ML15009A279
| ML15009A279 | |
| Person / Time | |
|---|---|
| Site: | 05000128 |
| Issue date: | 11/13/2014 |
| From: | McDeavitt S M Texas A&M Univ |
| To: | Wertz G A Document Control Desk, Office of Nuclear Reactor Regulation |
| References | |
| 2014-0069, TAC ME1584 | |
| Download: ML15009A279 (11) | |
Text
Jmj I TEXAS A&M ENGINEERING EXPERIMENT STATION NUCLEAR SCIENCE CENTER November 13, 2014 2014-0069 Document Control Desk ATTN: Geoffrey Wertz U.S. Nuclear Regulatory Commission Washington, DC 20555-0001
Subject:
Response to U.S. NRC Request for Additional Information, Review of The Fuel Pool Temperature on Fuel Temperature for the License Renewal for the Nuclear Science Center TRIGA Reactor (TAC No. ME 1584), from the Texas A&M University System, Texas Engineering Experiment Station, Nuclear Science Center Reactor (NSCR, License No. R-83, Docket 50-128)To Whom It May Concern: The Texas A&M University System, Texas Engineering Experiment Station (TEES), Nuclear Science Center (NSC, License No. R-83) operates a LEU, 1MW, TRIGA reactor under timely renewal. In December, 2003 the NSC submitted a Safety Analysis Report (SAR) as part of the license renewal process. In December, 2005 a conversion SAR (Chapter 18) was submitted resulting in an order to convert from the USNRC. In July 2009, the NSC submitted an updated SAR, dated June 2009, to the USNRC. This updated 2009 version of our SAR incorporated the information from the conversion SAR and the startup of the new LEU reactor core. On July 15, 2014 the U.S.NRC submitted a Request for Additional Information re: Review of The Fuel Pool Temperature on Fuel Temperature for the License Renewal for the Nuclear Science Center TRIGA Reactor (TAC No. ME 1584) as a part of the review process. Attached is our reply to this question.It is interesting to note that after the initial license application was made in 2009, a study was commissioned from General Atomics in 2010 to investigate the implications of a power uprate at NSC. Many of the issues regarding this RAI were carefully analyzed as part of that study and have been used as the basis for part of our response.
This response was reviewed internally by Dr. Karen Vierow, Associate Professor of the Texas A&M University Department of Nuclear Engineering.
We believe the contents of this response will satisfy the concerns raised by the reviewers.
If you have any questions, please contact Dr. Sean McDeavitt or Mr. Jan Vermaak at 979-845-7551.
TEL. 979.845.7551 I FAX 979.862.2667 nsc.tamu.edu 1095 Nuclear Science Rd. 1 3575 TAMU I College Station, TX 77843-3575 I declare under penalty of perjury that the foregoing is true and correct. Executed on November 13, 2014.Dr. Sean M. McDeavitt Interim Director, Nuclear Science Center Associate Professor of Nuclear Engineering Department of Nuclear Engineering Dwight Look College of Engineering Texas A&M University mcdeavitt@tamu.edu Xc: 2.1 1/Central File Duane Hardesty, USNRC Project Manager In response to the four items in the request for additional information (RAI) by letter dated July 15, 2014, the Nuclear Science Center (NSC) would like to withdraw the reference report "Evaluation of Pool Temperature on Fuel Temperature and MDNBR [Minimum Departure from Nucleate Boiling Ratio]"[1]
and replace it with "Safety Analysis Report for the Texas A&M University Nuclear Science Center Power Upgrade"[2], prepared for the Texas A&M Engineering Experiment Station by General Atomics under Research Service Agreement C13-00345 and dated.November 26, 2013.Justification The methodology used for the analysis presented in[1]involved the application of a single phase flow solver combined with a predictive correlation to determine the critical heat flux (Bernath correlation).
Upon further investigation, it was determined that the implementation was not subjected to extensive review to determine the validity of its application.
This is of particular concern when regarding the two phase flow regime existing during an approach to the Critical Heat Flux (CHF) which hints at the fundamental need for the flow solver to be able to calculate two phase flow. Stemming from this concern, the NSC will not continue to support the analysis done with this methodology and as a corrective measure will replace this supporting analysis with the safety analysis performed by General Atomics (GA) in November 2013 [2].The GA safety analysisis contained in a report [2] which was originally requested in order to support the operation of the NSC reactor at 1.25 MW, but with a nominal operating power of 1.0 MW. Additionally, the thermal-hydraulic analyses were performed for a pool inlet temperature of 60°C which was required to support a proposed limiting safety system setting (LSSS) on pool inlet temperature.
The GA safety analysis indicated an acceptable minimum DNBR of greater than 2.0 at a reactor power of 1.25 MW and a pool inlet temperature of 60°C.Summary of the relevant information in the GA safety analysis: In section 4.7 of[2], the analysis of the steady operation of the NSC reactor is presented.
In this safety analysis General Atomics has diverted from using their proprietary STAT code (as used in [3]) and replaced it with RELAP5/Mod 3.3 [4]for which a model description is provided[2].
The steady state analysis was performed for a pool inlet temperature of 60°C which differs that of [3] which was for a poolinlet temperature of 30°C. The 60°C inlet temperature, as well as an additional analysis at 1.25 MW, are parameters that were requested by the NSC as part of the service agreement to GA to allow for a more flexible basis for the reactor safety case (i.e. SCRAM limits). The following changes in operating parameters (between [2] and [3]) were observed:* Maximum clad outer temperature at 1.0 MW increased to 137.2°C from 135.9°C." Maximum fuel temperature at 1.0 MW increased to 526°C from 368.1°C.* Hottest channel outlet temperature at 1.0 MW increased to 101.0°C from 76.8°C.* Minimum DNBR increased to 2.53 at 60°C inlet temperature from 2.42 at 30°C inlet temperature.
The minimal increase in cladding outer temperature is expected due to the nature of the heat transfer properties at increased coolant temperature (i.e. dynamic viscosity decreases, thermal conductivity increases) as well as the higher thermal expansion coefficient of water at higher temperature (i.e. more buoyancy driven mass flow).The change in outlet temperature is directly related to the difference in inlet temperature (60°C vs 30'C) as well as minor differences in the total heat generation.
The only major change is that of the maximum fuel centerline temperature which increased by 158°C.This large step is counter-intuitive especially since the two models essentially should have the same solid conduction models, however, it was found that the step change resulted from the assumptions made for the gap between the cladding and the fuel pellets. In [3] it was assumed constant at 3 micrometers but in [2] it was specified as 12 micrometers at a reactor power of 1.0 MW which is 4 times larger. This increase in gap size dramatically increases the thermal resistance of the gap and translates to the associated increase in centerline temperature.
The change in minimum DNBR is largely due to the non-conventional method used in [3] to establish the DNBR, where the reactor power was increased to determine the point of unity in terms of DNBR.However, such a method unnecessarily penalizes the operational conditions and after internal investigations it was determined that, if [3] applied the conventional method for determining DNBR, the value would approximately be 3.32 at 30°C inlet temperature.
Also, the DNBR value at 60°C inlet temperature was also validated.
References
[1] General Atomics, "Evaluation of Pool Temperature on Fuel Temperature and MDNBR [Minimum Departure from Nucleate Boiling Ratio]", GA Project 30398, 30398R00001 Rev. 0, Nov. 2013.[2] "Safety Analysis Report for the Texas A&M University Nuclear Science Center Power Upgrade", General Atomics, Project 30398, November 2013.[3] "Safety Analysis Report -Texas A&M University System -Nuclear Science Center Reactor", Docket number 50-128, License number R-83, May 2011.[4] "RELAP5/Mod 3.3 Code Manual, Vols. 1-8", US Nuclear Regulatory Commission, NUREG/CR-5535, Rev. P3, March 2006.
Safety Analysis Report for the Texas A&M University Nuclear Science Center Power 30398R00001/0 Upgrade 4.7 Thermal Hydraulic Analysis -NSC LEU 4.7.1 Analysis of Steady State Operation
-NSC The following evaluation has been made for the TRIGA LEU 30/20 fuel system with four-rod configuration operating with cooling from natural colvection water flow through four-rod clusters of fuel. The steady state thermal-hydraulic performance of the NSC LEU TRIGA core was determined for operation at 1.0 and 1.25 MW with a water inlet temperature of 60 'C.The RELAP5 mod 3.3 computer code 4 was used to determine the natural convection flow rate, the coolant and clad axial temperature profiles, and the clad wall heat flux axial profile. The RELAP5 code was also used to determine the clad wall maxi -mum heat flux for departure from nucleate boiling. RELAP5 was also used to determine the fuel average and fuel maximum temperatures.
An average powered fuel rod and the maximum powered fuel rod were analyzed.
Four different flow geometries were used to analyze the maximum powered fuel rod so that the most limiting thermal hydraulic condition could be determined.
The arrangement of the four flow geometries around the maximum power fuel element (5D3) is depicted in Figure 4-2 1. Flow channel 221 is formed by the maximum rod 5D3, the transient rod 5D4, and fuel rods 5D I and 5D2. It has the smallest hydraulic diameter and smallest flow area. Flow channel 222 is formed by the maximum rod 5D3, the transient rod 5D4, and fuel rods 4D1 and 4D2. Flow channel 223 is formed by the maximum rod 5D3 and fuel rods 5E4, 4D2, and 4El. It has the largest hydraulic diameter.
Flow channel 224 is formed by the maximum rod 5D3 and fuel rods 5D2, SEl, and 5E4. The RELAP5 model shown in Figure 4-22 also includes flow channels for the instrumented fuel element (lFE). The fuel element with the maximum intra-rod power density is also shown in Figure 4-22 but is only used in the model for reactivity insertion transients.
41 Safety Analysis Report for the Texas A&M University Nuclear Science Center Power Upgrade 30398R00001/0 U ,tJ U All t~i'U ii;Figure 4-21: Flow Chnnnels That Include Maximum Fuel Element 5D3 42 Safety Analysis Report for the Texas A&M University Nuclear Science Center Power Upgrade 30398REJ0001/0 Figure 4-22: RELAP5 Model of Texas A&M NSC TRIGA 4.7.2 RELAPS Model Input -NSC The'RELAP5 analysis is a more modern and validated method than the method outlined in Section 4.6. The general reactor geometry and hydraulic data for the RELAP5 model are given in Table 4-16. The natural convection system for the NSC TRIGA is based on the four-rod cluster of fuel elements.
The representation used herein establishes multiple flow channels each typically bounded by four fuel elements.A RELAP5 steady state thermal hydraulic analysis was done for an average flow channel, an IFE channel and four flow channels containing the maximum powered fuel rod. The average flow channel has a flow area associated with the 90 fuel rods in the NSC LEU core. The IFE has the same hydraulic characteristics as the average flow channel except its flow area is for only one fuel rod. The hydraulic characteristics of the flow channels and power factorsof the associated fuel rods are presented in Table 4-17.The gap between the.fuel and cladding is an important parameter for predicting fuel temperatures.
Differences in thermal expansion between the fuel and cladding at operating temperatures reduce the initial manufactured gap and can result in gap closure at locations of 43 Safety Analysis Report for the Texas A&M University Nuclear Science Center Power 30398R00001/0 Upgrade high heat generation.
The average radial gap in the NSC LEU fuel is 2.15 rail (54.6 pul). The radial gap for the IFE in location 5E4 is 2.75 mil (69.9 pm). The radial gap for the maximum power fuel rod in location 5D3 is 2.66 mil (67.6 r.m). Table 4-18 presents the effective radial gap at various reactor operating power levels. For the maximum and IFE fuel rods, the gap size reduction is based on the peak axial power profile. For the average fuel element, the gap size reduction is based on an axial power factor of 1.0 so that the gap represents the average gap within the average fuel element.Table 4-16: RELAP5 Input for Reactor and Core Geometry and Hydraulic Data, NSC: Parameter Value Core and Reactor Geometry Unrteated grid plate length, in 2.00 Unheated lower refleclor length, in 3.95 Unheated top reflector length, in 4.06 Length from top of unheated length to top of pool surface, ft 23.685 Fuel element hetited length (in) 15 Fuel element diameter (in) 1.4126 Hydraulic Data Surface Roughness Grid plate K loss Ambient pressure at elevation 320 ft, psia Table 4-17: Hydraulic Flow Parameters and Power Factors, NSC..ara.eter .. :. ... !.' ig : I Channel Channel .Channel:.
Channel..22.1 222 223. 224.Flow area (in 2) 76.7196 0.852440 0.732495 0.929865 0.929581 0.735436 WVetted perimeter (in/element) 399.403 4.437814 4.494677 4.494677 4.437814 4.437814 Hydraulic diameter (in) 0.768342 0.768342 0.651877 0.827525 0.837873 0.662881 Fuel rod entrance K loss 0.236 0.236 0.174 0.281 0.281 0.176 Fuel rod exitX loss 1.289 1.289 0.952 1.534 1.533 0.960 Ma.ximum rod power factor 1.00 1.484 1.648 1.648 1.648 1.648 Power factor of other rods -- -- 1.584 1.432 1.400 1.493 Peak axial power factor 1.318 1.318 1.294 1.294 1.294 1.294 44 Safety Analysis Report for the Texas A&M University Nuclear Science Center Power Upgrade 30398RO0001/0 Table 4-18: Radial Gap Between Fuel and Cladding*Radial Cap in 1111 (11mr)Fuel Element Cold 0.5 o 1MW 1.0 MW .1.25 nV Average 2.15 (54.6) 1.55 (39.4) 0.95 (24.1) 0.77 (19.6)Instrumented Fuel Element 2.75 (69.9) 1.16 (29.5) 0.55 (14.0) 0.39 (9.91)Maximum Power Fuel Element 2.66 (67.6) 1.04 (26.4) 0.47 (I1.9) 0.33 (8.38)4.7.3 Steady State Analysis Results -NSC Tables 4-19 and 4-20 list the pertinent heat transfer and hydraulic steady state results for the 1.0 and 1.25 MW NSC TRIGA reactor, respectively.
Results are presented for an average chatnel and the four maximum powered flow channels (hot channels) at initial core conditions.
The average core temperature at 1.0 MW is 354.9 'C. It increases to 373.4 'C at 1.25 MW. The IFE located at 5E4 has a maximum temperature of 497.3 °C at 1.0 MW. It increases to 527.1 °C at 1.25 MW.Table 4-19: Steady State Results for NSC, 1.00 MW Channel Chann el Channel .'Channel*ý.lInitial core (90 Elemients)
Aerage 21222. :223 .224 Mass flow rate, kg/sec 0.0860 0.0875 0.0912 0.0984 0.0989 Exit coolant temperature, 'C 90.8 96.4 92.8 99.3 101.0 Coolant saturation temperature, °C 1 15.2 115.2 115.2 115.2 115.2 Maximum wall heat flux, WV/cm 2 34.09 55.16 55.16 55.16 55.16 Maximum flow velocity, cm/sec 16.20 19.26 15.76 17.18 21.78 Max. clad outer temperature, OC 131.9 136.7 137.2 137.2 136.6 Exit clad outer temperature, C 124.6 127.1 127.5 127.7 127.3 Peak fuel temperature, *C 478.7 525.3 525.7 525.6 525.2 Minimum DNBR -Bemath -- 2.58 3.05 2.92 2.53 Minimum DNBR- Groeneveld
'86 -- 4.79 4.88 4.75 4.75 Minimum DNBR -Groeneveld
'06 -- 4.24 4.04 3.94 4.18 45 Safety Analysis Report for the Texas A&M University Nuclear Science Center Power Upgrade 30398R00001/0 Table 4-20: Steady State Results for NSC, 1.25 MW Channel Channel Channel :Channel Initial Cot e (90 Elements)
'Average .1 222 223 224 Mass flow rate, kg/sec 0.0931 0.0998 0.1043 0.1170 0.1179 Exit coolant temperature, 0C 95.5 99.9 95.7 101.3 103.0 Coolant saturation temperature, °C 115.2 115.2 115.2 115.2 115.2 Maximum wall heat flux, \V/cnr 2 42.62 68.95 68.95 68.95 68.95 Maximum flow velocity, cmn/sec 17.61 22.19 18.22 20.72 26.38 Max. clad outer temperature, 0C 134.1 139.3 139.8 139.8 139.2 Exit clad outer temperature, 0C 126.4 129.0 129.2 129.5 129.1 Peak fuel temperature, 0C 516.5 561.6 562.0 562.0 561.5 Minimum DNBR -Bernath -- 2.02 2.39 2.31 2.01 Minimum DNBR -Groeneveld
'86 3.83 3.91 3.85 3.89 Minimum DNBR-Groeneveld
'06 -- 3.57 3.22 3.15 3.36 Tables 4-19 and 4-20 include results for the minimum departure from nucleate boiling (DNB)ratio using three correlations.
Historically, TRIGA reactors have used the Bernath correlation 5 to predict DNB. RELAP5 Mod 3.3 uses the 1986 Critical Heat Flux Lookup Table by Groeneveld 6.This Lookup Table was updated in 20067 to include more experimental data and to address concerns with DNB at low pressure.
The Bernath correlation is considered to be quite conservative and the Groeneveld 2006 correlation is considered to be more accurate.
All correlations have a minimum DNBR greater than the requirement of 2.0 given in NUREG-1537.
4.8 Thermal
Neutron Flux Valtes -NSC LEU Core 4.8.1 Thermal Neutron Flux Values in LEU Core, Calculated
[No change]4.8.2 Thermal Neutron Fltx Valtes in LEU Core, Measured[No change]46 Safety Analysis Report for the Texas A&M University Nuclear Science Center Power 30398R00001/0 Upgrade 1 Lawrence, R.D., "The DJF3D Nodal Neutronics Option for Two-and-Three-Dimensional Diffusion Theory Calculations in Hexagonal Geometry," Doc. No. ANL-83-1, Argonne National Laboratory, March 1983.2 Derstine, K.L., "DIF3D: A Code to Solve One-, Two-, and Three-Dimensional Finite Difference Diffusion Theory Problems," Doc. No. ANL-82-64 3 Los Alamos X-5 Monte Carlo Team, "MCNP -A General Monte Carlo N-Particle Transport Code, Version 5," LA-UR-03-1987, April 24, 2003.4, "RELAPS/Mod 3.3 Code Manual, Vols. 1- 8," US Nuclear Regulatory Commission, NUREG/CR-5535, Rev. P3, March 2006.s Bernath, L. B., "A Theory of Local Boiling Burnout and Its Application to Existing Data," Chem. Eng. Progress Symposium Series No. 30, Vol. 56, pp 95-116.6 Groeneveld, D. C., et al., "1986 AECL-UO Critical Heat Flux Lookup Table," Heat Transfer Engineering, Vol. 7, Nos.1-2, pp 48-62, 1986.7 Groeneveld, D. C., et al., "The 2006 CHF Look-up Table," Nuclear Engineering and Design 237, pp 1909-1922, 2007.47